Theoretical Formulation
4.2 Micro plasma arc welding
4.2.1 MPAW of Titanium alloy
In the present work, micro plasma arc welding is carried out on 500 μm thick Ti6Al4V, SS304 and low carbon steel individually. Figure 4.5(a) shows the weld beads of Ti6Al4V alloy obtained by bead on plate welding at a speed of 4.2 mm/s and current 11 A without the effect of extra shielding gas supply. The decolourization of the specimen is due to the formation of oxide film which leads to the degradation of mechanical properties. This is serious problem for welding titanium alloys. Several favorable activities have been followed to improve the joining system of titanium alloy using micro plasma arc welding process. In addition to proper shielding, the samples are cleaned using acetone solution to remove any oil attached and dried properly.
The use of intensive fixture is an essential requirement for producing a good weld joint. A TH-1698_11610311
suitable fixture was designed and constructed for the operation of welding procedure in open air to prevent both oxygen contamination and distortion of the workpiece. The fixture which is made of copper is shown in the Fig. 4.5(b).The gap left for welding is kept to a minimum only to allow the movement of the plasma torch. If the gap is large, the misalignment of the sheets at the joint occurs since the workpieces are very thin. The backing plate of the fixture is made of copper which ensures low angular distortions of the joint due to intensive heat transfer [Klimpel and Lisiecki, 2007]. Also the problem of oxidation is reduced due high rate of heat diffusion through copper. The flow rate of the shielding gas is kept higher side (~ 2 lpm) to protect the molten weld pool against oxidation. The flow rate is adjusted in such a way so that it protects the weld pool without blowing up of the molten metal. The welding electrode is made of Tungsten (W) with 2 wt. % Thorium (Th) of diameter 1.2 mm and the copper nozzle is of 1 mm diameter. The flow rates of plasma and shielding gas are selected as 6 lpm (litre per minute) and 0.6 lpm, respectively. The MPAW process is carried out at a range of current (7A-13A) and welding speed (2.33 - 6.67 mm/min) with a vertical torch position. Figure 4.5(c) demonstrates the welding cycle used in present case. The welding current is given to increase (upslope current) and decrease (downslope current) gradually during the welding procedure. This reduces the wear of the electrode and increases its lifetime. Both the upslope and downslope currents are taken as 4s.
Fig. 4.5 (a) Bead-on-plate welding at 11 A, 4.2 mm/s; (b) Copper fixture used in present investigation; (c) Welding cycle.
Autogenous welding is performed on square butt joint configuration on Ti6Al4V sheets of 100×50×0.5 mm at different power and welding speed based on the preliminary tests of bead on plate welding. The sheets to be joined are clamped tightly by a copper fixture so that there is minimum gap between two sheets. Edge preparation was done properly to keep the gap at the
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joint at a minimum level. The MPAW process parameters are presented in Table 4.4. Figure 4.6 (a) - (c) shows the weld joint obtained at different parameters. The surface seems bright silver that indicates there is no reaction with atmospheric gases during the cooling phase and during the actual welding. It is obvious that with increase in welding current the weld bead increases while an increase in welding speed decreases the weld bead. This is due to the fact that increase in current and decrease in speed leads to increase in heat input and subsequently results in bigger molten pool [Watanabe and Satoh, 1957]. Thus proper heat input is necessary to obtain a stable welding. From Fig. 4.5(b) and Fig. 4.6(b), it is observed that the weld bead decreases marginally with similar process parameters. This is due to presence of slight gap between two sheets (Fig.4.6b) for actual butt joint as compared to bead-on-plate welding. While identifying parametric envelops for a weld joint, three modes i.e. heating, stable and burnt through welding are identified to characterize the joint (Table 4.5).
Table 4.4 MPAW process parameters used in present investigation.
Welding Parameters Values
Welding current (A) 8 - 13
Welding speed (mm/s) 2.33, 2.75, 4.2, 5.62, 6.67
Copper Nozzle diameter (mm) 1.2
Electrode diameter(mm) 1.2
Plasma gas flow rate(lpm) 6
Nozzle to plate distance (mm) 2
Shielding gas flow rate (lpm) 0.4
Pre Flow (s) 4
Post Flow(s) 4
Torch Position Vertical
Welding current and speed are the most influential parameter because it affects bead shape, heat affected zone, the depth of penetration, and the amount of base metal melted. The total molten area increases with the increase in welding current. If the current is too high at a given welding speed, the fusion area will also be too high so that the resulting weld may tend to melt through the thickness. This over-welding increases weld shrinkage and causes greater
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distortion. If the current is too low, inadequate penetration or incomplete fusion may result. Too low current also leads to unstable arc, inadequate penetration, and overlapping. Figure 4.7 shows the feasible domain of welding speed and arc current for a successful weld joint. There is positive correlation between welding speed and welding current. It is found that at a speed of 2.33 mm/s and 8 A, a stable welding of the material occurs while at a speed of 6.67 mm/s and even at 11 A, no bead is formed and subsequently higher current is required to weld the material.
When the welding speed is increased, the proper welding current range shifts towards the higher- current side. The stable welding is obtained under the combination of weld current and speed.
Beyond a certain limit of the current, burn out of the material occurs. The selective welding speed and arc current indicates that there may be a process map that creates a domain of successful joints in plasma micro welding process. This graph can be used to determine the type of weld at different parameters and can further be used for predicting the types of welds at higher current and welding speed.
Fig. 4.6 Weld beads for butt joint obtained at: (a) 10 A, 4.2 mm/s, (b) 11 A, 4.2 mm/s, and (c) 10 A, 5.26 mm/s.
Table 4.5 Symbols used for categorization on type of welding.
Category Symbol Weld surface Cross-section
Heating ■
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Stable welding ●
Burnt through ▲
Fig. 4.7 Process map for micro plasma welding of Ti6Al4V butt joint configuration as a function of welding current and speed.
In arc welding, sufficient amount of energy has to be supplied to create the weld pool by fusion. The relative measure of the energy transferred per unit length of is heat input and is expressed as:
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(4.1)
where V, I, U represents the voltage, current and speed respectively. It is an important characteristic since it governs the cooling rate which may affect the mechanical properties and metallurgical structure of the weld and the HAZ. Fig.4.8 shows the influence of heat input per unit length (combined effect of welding power and welding speed) on weld dimensions. It is observed that at lower heat input weld bead increases more rapidly and the steepness decreases towards higher heat input. The trend indicates that the weld bead converges towards a fixed value with increase in heat input per unit length. Beyond certain heat input (~ 85 J/mm) there may not be any increment of weld width.
Fig. 4.8 Weld bead dimensions with heat input.
4.2.1.1 Macro and microstructural study
The detailed microstructural observation is conducted for each welded specimen using optical microscopy to determine the variation of grain size and presence of any weld defect. The microstructural analysis is preceded by fine polishing in emery paper followed by polishing in velvet cloth by suspension of alumina powder over it. Macro etchants differentiates major portions of specimen only like welded zone and normal surface. The macro etchant that is used
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for Ti6Al4V is HCL+H2O in 1:1 ratio. The micro etchant used is Kroll’s reagent i.e. distilled water, nitric acid, and hydrofluoric acid in the ratio of 46:3:1.
The strength and toughness of Ti6Al4V alloy are quite sensitive to microstructure.
Figures 4.9 and 4.10 show the macrographs for the welded Ti6Al4V sheets. The Fig.4.9 reveals the grain refinement at three zones corresponding to welding conditions of of 53.57 J/mmheat input per unit length. The boundary of the fusion zone is difficult to identify due to the narrow transformation kinetics for Ti6Al4V [Donaiche, 2000]. Macrograph demonstrates no sharp interface between FZ and HAZ suggesting epitaxial growth phenomenon in FZ from HAZ [Karimzadeh et al., 2005]. However, the shape of the fusion zone is regular and symmetric, and is clearly defined by semi-circular cross-sectional boundary. In all the cases, the fusion zone is elliptical in shape. It is found that the fusion zone size increases with increase in heat input. It is also reflected that full penetration and no unacceptable porosity and cracks are observed in the fusion (FZ) or heat-affected (HAZ) zone at any welds. This indicates that proper selection of welding conditions is implemented in all the cases. The depth/width ratio reveals a conduction mode welding for the investigated weld with elliptical shape of fusion zone. This is due to the fact that low heat input (~ 45-90 J/mm) of the plasma arc is not enough to produce a keyhole.
This conduction mode welding along with full and effective protection of the weld pool using argon shielding gas employed in the present case restricts high peak temperature of the system and reduces the reaction with air. Figure 4.10(d) shows a misaligned sample due to mis-fitting of the fixture. Excessive heat input during welding can cause considerable distortion, particularly in thin sheets. Thus very careful edge preparation and fitting of the sheets to be joined is required especially for thin sheets.
Fig. 4.9 Macrograph of weld cross-section at welding condition of 53.57 J/mm.
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After metallographic sample preparation, grains in a specific alloy are often analyzed via microscopy, where the size and distribution of these grains can demonstrate the quality of the sample. Grain size measurement is carried out by using line intercept method since the grains are not equiaxed. The
measurements are carried out using an optical microscope with digital photo capturing facilities
and analysis
software. In this method, five line of known length 'GL' is drawn and number of grain boundaries intercepted by the lines is counted
(designated as GN). The mean line average is given by:
(4.2)
Assuming reasonable shape of grains and obtaining a mean diameter of rotation, the average grain diameter is given by:
(4.3)
The weld fusion zone in titanium alloy is characterized by coarse, columnar, prior-beta grains that originate during solidification. In full penetration plasma welding, the columnar beta grains solidify inward from the base metal in a direction nearly parallel to the workpiece surface, ultimately impinging to form a vertical grain boundary at the weld centerline. The fusion zone beta grain size depends primarily on the weld energy input where a higher energy input promotes larger grain size. The weld joint mechanical properties, particularly ductility, can be degraded by a coarse prior beta grain [Karimzadeh et al., 2006]. Therefore, it is important to maintain fine grain structure by minimizing the weld energy input. In the heat affected zone, the grain size is significantly smaller than that of the fusion zone (Fig. 4.11a). The base metal exhibits an average Fig. 4.10 Cross-sectional macrograph at welding condition of: (a) 48.7 J/mm; (b) 65.5 J/mm; (c) 72.72 J/mm; (d) 85.5 J/mm.
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grain size of ~ 15 μm. From Fig. 4.11(a) it can be observed that very large prior-β grains in the FZ with an average size of 214 μm exist. It is also observed that a significant grain growth in the HAZ where the average prior-β grain size is~ 67 μm. The evolution of the grain size of the prior- β phase in the HAZ and the FZ with respect to heat input is represented in Fig. 4.11(c). The change in heat input causes some prior-β grain refinement in the FZ. Indeed, the average of prior-β grain size increases from 151 µm to approximately 297 µm when the heat input increases from 44.4 J/mm to 96.6 J/mm. It is also noteworthy that minor refinement of prior-β grain is observed in the HAZ due to change in heat input. In this region, the average prior-β grain sizes that are corresponding to the lowest (~ 44.5 J/mm) and the highest (~ 96.6J/mm) heat input are of 54 µm and 74 µm, respectively. However, the refinement of grains in HAZ zone is not as significant as in the fusion zone. These results demonstrate that the heat input strongly affects grain growth and the large grain size can be avoided by using smaller energy input whenever possible. In general, large temperature variation takes place within a small length along its width.
The molten zone is subjected to a very high temperature which is confined within a small area.
This promotes grain coarsening at the fusion zone. Also due to the low thickness of the plate, the heat flow is limited along the width direction. Thus, wide variation of grain size occurs along the width. However, increasing the speed or reducing the heat input per unit length, can reduce the difference of grain size in the three zones as is reflected in Fig. 4.11(b).
Fig. 4.11 (a) – (b) Macrographs showing different zones for Ti6Al4V weldment at heat input of 65.5 J/mm; (c) Prior-β grain size in the FZ and HAZ for the different welding conditions.
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In welding of Ti6Al4V, 𝛼→β phase transformation during rapid heating and the decomposition of the β→𝛼 phase during cooling greatly depends on heating or cooling rate. If the cooling rate is higher than the critical cooling rate (~ 623K/s) of Ti6Al4V, 𝛼' martensite is formed (diffusionless transformation). At slower cooling rate, diffusion controlled nucleation takes place resulting in the growth of secondary lamellae [Fan et al., 2005]. Some β phase is always retained irrespective of the cooling rate. Due to the high thermal conductivity of copper, the cooling rate of the material is considerably increased. However, the typical cooling rate is
∼100 K/s for PAW process which is approximately increased to 700 K/s using copper as a fixture which greatly affects the formation of microstructure and is sufficient to form a martensitic structure. The microstructure observed in Fig. 4.12 reveals that all these samples exhibit very similar microstructures in the fusion zone consisting of fine acicular martensitic 𝛼'structure within the β grains. This microstructure forms when the structure is quenched from the β phase above the beta transus temperature (1253K), at a cooling rate higher than the critical cooling rate [Ahmed and Rack, 1998]. Thus, it may be concluded that there is not much difference in the cooling rate in the investigated region. Although each location contains 𝛼' martensite, the amount, size and distribution differ at each location. The 𝛼' phase is a martensitic phase with an HCP crystal structure and similar lattice parameters to the 𝛼 phase. The dark etching phase appears within the martensitic structure is retained β phase. However, the microstructure in different zones is indirectly dependent on heat input.
Figure 4.12(a) shows the microstructure in the fusion zone at heat input of 48.7 J/mm.
At this magnification, the microstructure appears to be composed entirely of martensite oriented in different direction in prior beta, where the length of many of the needles exceeds 100 µm. This microstructure has very less retained β, and appears to form entirely through diffusionless martensitic transformation. Figure 4.12(b) represents the microstructure of fusion zone at heat input of 53.6 J/mm where length of the 𝛼' martensite increases and the distributions becomes finer that exhibits relatively better strength [Ahmed and Rack, 1998].It is observed that martensitic length decreases with further decrease in heat input (Fig. 4.12c) corresponding to heat input of 65.5 J/mm. At a higher heat input (Fig. 4.12d), it shows completely different microstructure. The formation of a second morphology i.e. α-β lamellar structure with α-phase lamellae in a β-phase is observed in the fusion zone. The 𝛼 lamellae is arranged in a Widmanstatten/basket weave structure also called Thomson structures with different sizes and
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orientations, and forms alpha platelet colonies within the columnar grains [Rafi et al., 2003].
This may be presumably due to moderate cooling rate slightly lower than critical cooling rate because of the high heat input [Elmer et al., 2003]. The fusion zone at high heat input shows mixed mode of transformation which consists of initial mode of transformation and subsequent martensitic transformation. Such structure has lower strength than the 𝛼' martensitic structure [Ahmed and Rack, 1998]. Thus, it may be concluded that the percentage of martensitic structure decreases with increase in heat input. The lowest heat input which has the highest cooling rate has the highest martensitic fraction. Only 𝛼' is present in the fusion zone formed by the lowest heat input while both 𝛼' and transformed 𝛼+β are formed in the fusion zone obtained from highest heat input [Elmer et al., 2003]. From Fig. 4.12(a) and Fig. 4.12(b) it can be concluded that the average martensite length increases from 50 µm to 100 µm, however the density becomes coarser. Again from Fig. 4.12(c) and Fig. 4.12(d) shows that the average martensite length reduces to 30 µm.
Fig. 4.12 Microstructure of fusion zone for Ti6Al4V weldment at welding conditions of: (a) 44.96 J/mm (12A, 6.67 mm/s); (b) 53.6 J/mm (9 A and 4.2 mm/s); (c) 65.5 J/mm (11 A and 4.2 mm/s); (d) 85.83 J/mm (8 A, 2.33mm/s).
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The peak temperature and cooling rate decreases away from the weld centerline. The heat affected zone (HAZ) experienced significant phase transformations during the welding process due to different peak temperature, and heating and cooling rates. It is observed from Fig. 4.13(a) and Fig. 4.13(b) that far away from the weld center line, the blocky secondary phase is produced under a lower cooling rate along with some primary 𝛼 and few traces of retained β . However, Fig. 4.13(c) reveals that the martensitic 𝛼' is substantially formed due to high cooling rate at the near HAZ zone. It is also observed that the content of acicular 𝛼' is decreased significantly from the region near the fusion zone to region near the base material. This is because that the high temperature gradient exists in the narrow HAZ due to low heat input of the process (48.7 J/mm).
Fig. 4.13(b) demonstrates that the primary phase exists in the middle of the HAZ that contains secondary Widmanstatten 𝛼, having morphology of a blocky appearance with a heavily dislocated internal substructure and with a small amount of primary 𝛼. The appearance of the original 𝛼 phase and β phase in the HAZ signifies partially transformed region [Elmer et al., 2003] i.e. the region may not be heated to the beta transus temperature (1253K). Fig.4.14 highlights the microstructure of HAZ zone corresponding to the highest heat input of 96.6 J/mm.
The HAZ for this welding condition is entirely different from low heat input (~ 48.7 J/mm). This is because of high heat input causing slower cooling rate resulting in diffusion controlled transformation. Figure 4.14(a) shows the near-HAZ region which also has a microstructure similar to the fusion zone, consisting almost entirely of the transformed grains with a lamellar structure. With further away from the fusion zone as shown in Fig. 4.14(b), an increase in the volume fraction of this morphology is observed due to the reduction in cooling rate resulting in the further growth of the grain boundary Widmanstatten 𝛼 plates into the center of grains and formation of the classical ‘basketweave’ Widmanstatten morphology. In the far-HAZ region (Fig. 14.14c), with increasing distance from the FZ boundary, the fraction of the transformed β with a lamellar structure decreases whereas the primary equiaxed 𝛼 fraction increases.
Fig. 4.13 Microstructure of the HAZ for heat input of 48.7 J/mm.